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Thermochemical analysis of ACFB-based gas and power cogeneration
Pergamon
Enrr~’ Vol. 20, No. 12, pp. 1X-1283. 1995 Copyright 0 1995 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0360.5442/95 $9.50 + 0.00
0360-5442(95)00063-l
THERMOCHEMICAL ANALYSIS OF ACFB-BASED POWER COGENERATION S. V. MAKARYTCHEV,?
GAS AND
M. X. FANG, X. T. LI, Z. Y. LUO, and K. F. CEN
Institute of Thermal Power Engineering, Zhejiang University.
3 10027 Hangzhou. People’s Republic of
China (Received 23 December
1994; received for publication 7 June 1995)
Abstract-A first- and second-law analysis is presented for a process developed for simultaneous generation of a fuel gas and electric power (gas and power cogeneration) based on atmospheric circulating-fluidized-bed (ACFB) combustion of coal. The mathematical model has a zone structure, multi-species equilibrium calculations for applicable zone conditions at high temperatures (50 gas-phase and seven solid-phase chemical species) and the concept of freezing of the gas composition at low temperatures. Our analysis shows that the process utilizes coal in a simple, effective and environmentally clean manner. The first- and second-law efficiencies of the process are, respectively, 35.0 and 27.6% for gas generation, 15.4 and 14.6% for power generation, 50.4 and 42.2% overall. The heating value of the gas is 11 MJ/Nm3 (medium). Desulphurization is achieved by using CaS-based sulphur capture during limestone addition to the gasifier bed. Results are compared with data from a 150 kg of coal/h experimental plant.
INTRODUCTION of replacing liquid and gaseous fuels derived from petroleum and NG by synthetic fuels from coal is generally recognized. The technological basis for utilizing coal in combustion and converting it to fuel gas has been created. I,* However, in view of the wide technological gap between industrialized and developing countries, it is unrealistic to expect worldwide use of these highly effective but expensive technologies in the near future. In some rapidly developing countries (e.g., China), there is a need for new gas- and power-generation techniques adapted to the existing technological level. Recently, much effort has been expended in China to develop circulating fluidized bed (CFB) technology to utilize vast national coal resources in an inexpensive, effective and environmentally clean manner. CFB-based power generation in China is now reaching the status of a commercial technology.4 Based on experience in CFB-related research and development, the Institute of Thermal Power Engineering of Zhejiang University started a project aimed at development of a process for simultaneous generation of a fuel gas and electric power (gas and power cogeneration). The first stage is based on ACFB combustion of coal leading later to a pressurized operation. A description of the 150 kg,,,,/h experimental ACFB-based gas/steam generator (the key component of the gas and power cogeneration plant) has been given by Cen et al,’ and a process analysis was carried out by Makarytchev and Cen.‘j A 14 t_,,/h plant is currently under construction in Yangzhong county, Jiangsu province, China, and is scheduled to start operation by the end of 1995. In this paper, we present an analysis of an ACFB-based gas and power cogeneration plant for commercial applications. Because of the heterogeneous nature of the process, the analysis is based on a thermochemical approach, i.e. changes of the thermodynamic quantities within the system (enthalpy, entropy, exergy, etc.) are derived from transformations of chemical compositions and the thermodynamic state of the multicomponent energy carriers, which include fuel and flue gases, solids, and water or steam. The emphasis is on gas parameters and plant efficiencies.
The desirability
PROCESS DESCRIPTION Figure
gas/steam
1 shows the general configuration of a gas and power cogeneration plant. It comprises a generator and a steam turbine cycle. The steam cycle of the plant is a traditional one; it
tvisiting research professor (formerly Moscow University, Russia). To whom all correspondence should be addressed. Present address: Department of Chemical Engineering, University of Sydney, N.S.W. 2006, Australia. 1271
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S. V. Makarytchev
et al
Fuel gas Pump
A
I Steam Solids Flue gas to stack N Feed water
Char 1
1
ACFB combustor
Air
Fig. 1. Configuration
ST
Boiled superheater
Steam
of the gas- and power-cogeneration
plant.
consists of a steam turbine, a condenser, a feed pump and a steam rising circuit (economizer, boiler and superheater). A dual-fluidized-bed gas/steam generator (Fig. 2) has two interconnected reactors: an FB gasifierpyrolyzer and an ACFB char combustor. The gasifier bed is fluidized by the mixture of steam, supplied by an internal steam generator, with a recycled part of a fuel gas. The combustor bed is fluidized by air from a compressor. Crushed coal is supplied to the gasifier along with limestone for coal sulphur retention. In the gasifier, coal is heated, pyrolyzes, and reacts with the steam. Hot fuel gas is cooled in an economizer section of a convective pass and, after filtration and dehydration, it is accumulated in storage for future utilization. The char formed during coal pyrolysis is moved to the char combustor by the circulating inert solids and is burned there. The upper part of the combustor contains a superheater for a steam-rising circuit. After the superheater, hot flue gas is separated from solids by a cyclone-type separator and is directed to the convective pass, which contains an internal steam generator and a boiler. The hot solids from the separator recycle to the gasifier and provide heat needed for coal pyrolysis and gasification. The thermal balance of the dual-fluidized-bed combustor-gasifier system may be described as follows. An overall plant thermal input with coal HVcoal is split by the gasifier into two parts: a combustor thermal input fi, (char) and a gasifier thermal input E,, i.e. HVcoal= *” + @” .
(1)
The ratio of fi, to fi, is closely related to the gas-generation process and will be considered below (see Table 3). The heat released in the combustor is used for heating circulating solids AQs, raising steam in the superheater Ae,H, and increasing the flue-gas enthalpy AHfg, viz., H$ = AQs + Ae,n + AHfg .
(2)
I213
ACFB-based gas and power cogeneration ACFB combustor
Gas/solid seperator
Internal steam generator
Flue gas
Fig. 2. Layout of the gas/steam generator.
The gasifier converts its thermal input and the heat delivered by the circulating solids to the fuel-gas heating value HV,,, and a fuel-gas enthalpy AH,,,, i.e. en + AQs = HV,,, -I- AHgaS.
(3)
The combustor-gasifier temperature difference AltG and the solids circulation rate must be sufficiently great to provide the required heat transfer from the combustor to the gasifier. The optimal temperature for the CFB combustor is determined by the level of sulphur retention and an acceptable NO, emission level; it is known to be 850-950°C3 The lower limit of the gasifier temperature is dictated by the rates of pyrolysis and gasification reactions.7 Once the reactor temperatures are chosen, a thermal balance is achieved by adjustment of the solids-to-coal mass-flow ratio, viz.,
kodkgcoa, = AQs/(cs AF=) ,
(4)
where cs is the specific heat of the circulating solids. The present design is based on temperatures of 950°C for the combustor and 8OO’C for the gasifier. For these temperatures and 0.3 mm silica sand with cs = 0.81 kJ/kg K, Eq. (4) yields a lower bound for the solids-to-coal mass-flow ratio of 7.4 kgsorids/kgcoar. The actual ratio should be somewhat higher because of inevitable heat losses. MATHEMATICAL MODEL AND ASSUMPTIONS
The analysis is based on a thermochemical approach, i.e. changes of the chemical compositions and thermodynamic states of the multicomponent energy carriers are used for calculations of the thermodynamic values. The first-law analysis refers to enthalpy transfer within the gas/steam generator and the steam cycle. The second-law analysis relates to exergy transformations.”
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S. V. Makarytchev et al
Gas/steam generator
The mathematical model of the gas/steam generator has a zone structure and requires equilibrium calculations for the applicable conditions. In the gasifier and combustor, the chemical compositions of the hot gases are assumed to be in equilibrium at the local temperatures. During the temperature decrease in the heat exchangers of the convective pass, the gas composition changes but remains at equilibrium at the local temperature. At a certain section of the convective pass, the reaction rates become insufficient to achieve equilibrium and the gas composition then remains unchanged downstream. Comparison of the characteristic times for reactions with residence times of the gas shows that chemical composition freezing occurs close to the combustor and gasifier exits. Therefore, the gas composition was treated as frozen starting from the boiler and the economizer sections of the convective pass. The stack temperatures for both fuel and flue gas were taken to be the same at 200°C. The assumed environment for the second-law analysis was as follows: temperature = 25”C, pressure = 1 atm, mole fractions of the gas-phase components are 0.7561 for N1, 0.2034 for 02, 0.0312 for H,O(g), and 0.0003 for CO*. The condensed phase was assumed to be composed of H,O(l), CaC03(s), and CaSO, x 2H,O(s). Calculations for the gas/steam generator were carried out by using computer code for determination of the gas- and condensed-phase species concentrations and thermodynamic functions in multicomponent systems’s” adapted to fossil-fuel applications. For the coal/steam/limestone/air system of concern, the 57 chemical species included were the following: gas-phase Hz, H, CH, CH2, CH3, CH4, C2H, C2H2, C2H3, GH4, C2H5, CzHs, NZ, N, NH, NHZ, NH3, N2H2, CN, COS, NCO, HCN, 02, 0,03, OH, HOZ, Hz02, H20, CO, CO*, HCO, COOH, CH*O, CH30, NO, NO,, NzO, NZ03, N204, NzOs, HNO, SZ, S, SO, SOZ, S03, HS, H,S, Ar; solid-phase C, Ca, CaO, CaS, CaCO,, CaSO, and CaSO,. Data for the species were taken from thermodynamic tables of Glushko. I’Details on chemical equilibrium calculations may be found elsewhere.12*i3 Steam cycle
Assumptions concerning steam-cycle parameters, which are close to design values, are shown in Table 1. Thermodynamic data for steam and water were taken from the steam tables of Keenan et al as given by Burghardt.14
FUEL-GAS GENERATION
Reactions in the gasifier and combustor
When the crushed coal particles are entrained in the steam-blown gasifier, the following reactions occur in succession: coal drying and pyrolysiscoal(s) - msH*O(g ) + Ash(s) + fcC( s) + vmC,H,,OJ,S,( moisture char volatiles
Table 1. Assumptions concerning the steam cycle. Parameter Steam condition pressure ( MPa) temperature (“C) Condenser pressure (kPa) Cooling water temperature (“C) entrance exit Steam turbine isentropic efficiency Feed pump efficiency Heat exchangers efficiency
Value
3.81 450 8.5 25.0 32.0 0.85 0.70 0.80
g);
(5)
ACFB-based
Table 2. Determination
gas and power cogeneration
of the extent of coal conversion Calculated
Experimental composition (wt%)
Gas component
co
I unreacted 26.46 3.49 15.57 0.01 1.88
36.19 5.63 13.56 2.61 2.69 I .69
H? CH., GH, N? I-M other
gas-phase
composition
(wt%) III reacted VM + IO%FC
36.88 8.76 6.25 -
47.13 7.38 10.23 -
1.72
1.60 I .46 16.67 14.26
1.62
0.77 49.66 0.70
31.34 0.23
in the gasilier.
II reacted only VM
1.46
co2
Hz0
VM + steam
1275
19.21 24.93 0.63
1.27
reactions of volatiles with steam-
C,.H,O,N,S, heterogeneous
+ (c-fc-o)HIO
carbon-steam
-
(c-fc)CO
+ (c-fc+h/2-o-s/2)H,
+ (n/2)N,
+ sH,S;
(6)
reactions (char gasification)-
C(s) + I-VW - CO(g) + Hz(g) CO(g) + H,O(g) us> + CO*(g) -
(7)
9
CO, + H,(g) 9
(8)
2CWg) .
(9)
In the preceding equations, C(s) is solid carbon bound in the char along with ash. The coefficients ms, fi, vm, and c, h, o, n, s in Eqs. (5) and (6) denote, respectively, the mole fractions of moisture MS, fixed carbon FC, volatile matter VM, and individual elements C, H, 0, N, and S in the raw coal. These coefficients are determined by the coal composition (see Table 3). The volatiles are composed of gaseous CH4, C2H6, CO, CO*, HZ, H20, H2S, NH,, and some other compounds. Because of the short residence time, the slow heterogeneous reactions (7)-(9) are not completed in the gasifier. Along with the circulating inert solids, the unreacted char moves to the combustor where heterogeneous char combustion occurs as follows: Table 3. Fuel characterization Reactor Parameter
Plant
Gasifier
Combustor
Fuel type Fuel composition
raw coal
VM+MS
FC + Ash
50.56 25. I 1 2.71 21.62 63.39 3.88 6.51 1.13 0.76 24.58 1.00 24.58 -
-
70.05
(wt%) FC VM MS Ash C H 0 N S Heating value (MJ/kg,.,)
Mass input (kg,,&,,,) Thermal input (MJ/kg,,,,,) Theor. air (kg,,$kg,,,) (kg,i,lkg,,,) Theor. steam (kg.,,.,/kg,.,) (kg,,,.,/kg&
90.26 9.74 46.12 13.95 23.40 4.06 2.73 28.75 0.278 8.00 0.43 0.12
-
29.95 70.05 -
22.91 0.722 16.58 8.06 5.82 -
S. V. Makqtchev
1276
C(s)
+ O*(g)
C(s) + (wwg) CO(g) + (1~2Mg)
et al
-
CO,(g)
(10)
9
+ CO(g) 9
(11)
-
(12)
a(g)
*
Extent of coal-to-fuel-gas conversion The extent of this conversion and the chemical composition of the fuel gas depend on the gasifier temperature, pressure and the residence times of reactants. For heterogeneous gas-solid reactions, additional important factors are coal reactivity, heating rate, reactive surfaces, and surface temperatures. The design operating temperature of the gasifier (NWC) and the residence time of the coal particles (approximately, 5 min) are sufficient for complete pyrolysis [reaction (5)] but not for completion of the char-gasification reactions (7)-(9), which are slow at 8OO’C.’ To determine the extent of reaction in the gasifier, the composition of the fuel gas generated in a 150 kg,,,,/h experimental gas/steam generato? was compared with equilibrium calculations. Table 2 shows the experimentally observed gas composition and the composition calculated for the three suc-
cds rslio 0
t
i 10’
0
0.4
0.8
1.2
1.6
2.0
Gasifier steam ratio 1.0
(b)
3 E g
0.5
2
Cold 9~ efficisacy
.i 3
Gasifies: MOT! Combttstor 95OT. air rstio 1.2
0
-
I
I
I
I
0.4
0.8
1.2
1.6
20
2.0°
Gasifier steam ratio Fig. 3. Effect of gasifier steam on fuel-gas composition (a) and coal-to-gas conversion parameters(b).
1277
ACFB-based gas and power cogeneration
cessive gasification steps: (I) only pyrolysis reaction (5) is completed (unreacted VM and steam); (II) the gas-phase reaction (6) of the volatiles with steam is completed (only VM is reacted); (III) the gassolid char-gasification reactions (7)-(9) are partially completed with 10% of the total char mass gasified (VM and lO%FC are reacted). The assumption for step III was used previously by Nag et alI5 for analysis of a pressurized, air-blown, Lurgi-type FB gasifier. Table 2 indicates that for atmosphericpressure, steam-blown gasification at 800°C, the observed concentrations approach the calculated values for step II. This means that the gas-phase reactions are close to completion while the extent of char gasification is negligible. The gasifier acts as a pyrolyzer-carbonizer. In view of these results, we have based the plant analysis on the following assumption: the fuel gas at the gasifier exit is an equilibrium volatiles-steam mixture at 800°C after completion of reactions (5) and (6) and all of the char formed during pyrolysis is burned in the combustor. For these assumptions, fuel characteristics for the gasifier and combustor are given in Table 3. Theoretical requirements for gasifier steam and combustor air are determined according to reactions (6) and (lO)-( 12), respectively. Fuel-gas
parameters
The main parameters determining gasifier performance are the fuel-gas composition, the contents of carbon in the fuel gas (kg/Nm3), the gas heating value HV,, (kJ/Nm3), and the gas yield YgaS (Nm’/kg,,,,). The main dimensionless parameters are the carbon conversion (ac) and the cold fuelgas efficiency niaS(first-law value); these are defined as follows: ffc = carbon-in-fuel-gas x Y&carbon-in-coal ,
(13) (14)
rltas = HV,,S Y&?SlHV,,,. Using the data of Table 3, carbon conversion in the gasifier is approximated as oc = (C-FC)/C = (63.39-50.56)/63.39 = 0.20 .
(15)
The fuel-gas composition and cold fuel-gas efficiency as functions of the gasifier steam ratio (actual steam/stoichiometric steam) are shown in Figs. 3(a) and 3(b). With an increase of the steam ratio, the concentrations of the combustible species CO and CH, in the fuel gas decrease, while those of the oxidation products HZ0 and CO* rise [Fig. 3(a)]. This explains a decrease in the gas heating value from 16 MJ/Nm3 for zero steam ratio (dry coal pyrolysis) to 9 MJ/Nm3 for a steam ratio of 2.0 [Fig. 3(b)]. At the same time, the gas yield increases from 0.5 to Sulphur capture --------ce __ ___--_-
106
,
.@
1.0
#*
s c ,a 0.4 g ?
-
‘J
stoam rrlio 1.O
0
1
2
3
4
Co/S molar trtio Fig. 4. Effect of limestone on H2S content in the fuel gas.
5
1278
S. V. Makarytchev et
al
1.ONm-‘/kg,,,,. As the result, & remains almost unchanged at 35% for the steam-ratio range O-2.0. At the design steam ratio of 1.O,IN,,, = 10.9 MJ/Nm3 (2600 kcal/Nm”), which corresponds to medium heating-value gases and agrees well with the value 10-12 MJ/Nm3 for the experimentally generated gas.5 Desulphurization
During pyrolysis, volatile sulphur compounds are produced as H,S and a small amount of COS. Even for low-sulphur coal, the amount of H2S in the fuel gas is about 2% [Fig. 3(a)], which poses an environmental issue. Desulphurization of the fuel gas is achieved by adding a Ca-based sorbent (in the present design, limestone) directly to the gasifier bed. At the gasifier temperature, limestone calcifies, i.e.
0.8
0.2
0
0.4
0.8 Gasifier
1.2
1.6
2.0°
steam ratio
(b) 1.0
0.8
0.6
u.6
1.0
1.2
I .4
1.6”
Combustor air ratio Fig. 5. Effects of gasifier steam (a) and combustor air (b) on the thermal balance of the gas/steam generator.
ACFB-based
gas and power cogeneration
1279
CaO( s) + COz( g) .
(16)
CaCO,( s) CaO then reacts with H2S as follows: H&g)
+ CaO( s) -
H,S(g) + (3/2)0,(g)
CaS(s) + H,O(g)
,
(17)
SO,(g) + H,O(g)
7
(18)
-
SO,(g) + CaO( s) + 3CO( g) -
CaS( s) + 3C02( g)
(19)
In an oxygen-deficient gasifier atmosphere, the amount of SO2 is negligible and sulphur is primarily converted to solid CaS. Sorbent particles containing CaS are transferred by the circulating solids to the combustor where CaS is oxidized to solid CaSO, by the process
Condensate 4.1
I
Combustor
Economizer
v////////-//m
Fuel gas
I
Boiler + superbeater
I
Steam turbine
-
I
I
Power 15.4 Fig. 6. Plant heat-flow diagram.
Interoal
Stackgas and loss
S. V. Makarytchev et al
CaS( s) + 202(g) - CaSO,( s)
(20)
and are then disposed of with the ash. Figure 4 shows the calculated amounts of H$, CaO, CaS in the gasifier and the fractional sulphur capture as CaS vs Ca/S molar ratio of an added sorbent. By increasing the amount of sorbent, the H,S concentration may be reduced to meet environmental constraints. However, increased sorbent addition results in larger amounts of solid waste and also leads to an increase in NO, emissions.‘6*‘7The optimal value for the Ca/S ratio is in the range 2-2.5. Details of the CaS-based desulphurization in a fluidizedbed reactor are considered elsewhere.13 RESULTS OF THE FIRST-LAW
ANALYSIS
Thermal petiormance of the gas/steam generator
The gas/steam generator converts the plant thermal input HVcoalto the fuel-gas heating value HV,,, in the gasifier and to steam heat in the economizer AHzc, boiler AZ,, and superheater AHzH.The thermal balance equation for the gas/steam generator is HV,,,, = HV,,, + W?
+ A@, + A@? + Q,oss9
(21)
where Qlossis the total heat loss in the heat exchangers and with the stack gases. Figures 5(a) and 5(b) show the normalized on HV,,, terms of Eq. (21) as functions of a gasifier steam ratio [Fig. 5(a)] and a combustor air ratio [Fig. 5(b)]. The relation between the terms of Eq. (21) is stable over a wide range of operating parameters. The increase of Qlossat the air ratio less than unity is caused by the rise of the content of combustible matter (mainly CO) in the flue gas. In a design operating regime (gasifier: 8OO”C,steam ratio 1.0; combustor: 95O”C, air ratio 1.2), 35% of the HVcoa, is converted to the HV,,. 45% is used for rising steam (2, 17 and 26% in economizer, boiler and superheater, respectively), and 20% is lost in the heat exchangers and with the stack gases. The first-law thermal efficiency of the gas/steam generator T& is the sum of the gas-generation (vi,,) and steam-generation efficiency ($), viz.,
it equals 0.35 + 0.45 = 0.80. 80 -
- 80 Comburtor: 9SO°C Guifier: 800°C, rtoam mtio 1.0
60-
-60
2 6 I 3 s u 1
Ovomll officioncy
40 -
-
Gas generation
Powor gonoration
0 0.6
I
I
I
I
0.8
1.0
1.2
1.4
Comburtor
air ratio
Fig. 7. The effect of combustor air on the plant thermal efficiency.
-0 1.6
40
ACFB-based
1281
gas and power cogeneration
Gasifier 1
I
Condc
r
Solids
oadenser 1.7 Power 14.6
Fuel gas
Fig. 8. Plant exergy-flow
Power generation
diagram.
and overall plant eficiency
Figure 6 illustrates the energy conversion process within the plant. The numbers in Fig. 6 correspond to the enthalpy values of the different energy carriers in the design operating regime. Regarding the steam cycle efficiency, only 31% of the superheated steam heat ( 15.4% of the HV,,,,) is converted to the steam turbine shaft work AWsTwhile the rest is withdrawn by the cooling water in the condenser. The first-law-based indexes of plant performance are the cold fuel-gas efficiency qk,, [Eq. (14)], power-generation efficiency
and overall plant efficiency I I I qo = qgas + qpower ’
(24)
The values of these efficiencies are almost independent of the gasifier steam ratio. Their dependence
S. V.
1282
Makarytchev et al
on combustor air ratio is shown in Fig. 7. In the design regime, the first-law efficiencies are 0.350 for gas generation, 0.154 for power generation, and 0.504 overall. RESULTS OF THE SECOND-LAW ANALYSIS
Exergy conversion,
losses and irreversibilities
Figure 8 shows the process of the exergy conversion in the gas and power cogeneration plant under the same conditions as for Fig. 6. Black triangles in Fig. 8 indicate the destruction of the exergy due to irreversibilities in the plant components. Table 4 gives the summary of the exergy analysis. Results show that 27.6% of the coal chemical exergy Scoalconverts to chemical exergy of the fuel gas ggas and 14.6% to steam turbine work AWsT. The exergy losses through the effluent streams are insignificant: 3.4% with the stack gases and 0.3% with the cooling water. This is in contrast with the corresponding first-law-based data: 5.8 and 24.0% of the HV,,,, respectively. The main source of the exergy loss (50% of the ‘%,,,,) are the irreversibilities within the gas/steam generator. Two-thirds of this amount is the result of imperfect heat transfer in the heat exchangers, and the rest is an internal irreversibility of the gasification and combustion processes. Second-law eficiency for gas and power cogeneration
Using data of Table 4, the second-law efficiencies may be evaluated as follows: qi= = C%gaslc&coa, = 0.276 , II
qPwer = AWsTIZ,,, = 0.146 I$
=
qias
+
q;,,,
=
(25) ,
(26)
0.422 .
(27)
Because of irreversibilities, these efficiencies are smaller than the corresponding first-law efficiencies. To improve process efficiency, there must be a decrease in the irreversibility of gas-to-steam heat transfer in the boiler and the superheater (32% in the total exergy balance). This may be accomplished by avoiding excessive heat dissipation into the environment by using better isolation materials and by appropriate matching of the temperatures of heating and heated streams. The irreversibilities in gasifier (7.5%) and combustor (9.8%) are inevitable features of the related chemical transformations. They can
Table 4. Summary of the exergy analysis. Parameter Exergy input with coal
26007
100.0
Exergy conversion fuel gas chemical exergy turbine work output total conversion
7174 3797 10971
27.6 14.6 42.2
892 77 969
3.4 0.3 3.7
1948 2524 5447 2875 252 669 338 44 14067
7.5 9.8 20.9 10.9 0.9 2.6 I.3 0.2 54. I
Exergy losses with the stack gases with the cooling water total losses Irreversibilities gasifier + economizer combustor superheater boiler internal steam generator turbine condenser feed pump total irreversibil.
ACFB-based gas and power cogeneration
1283
be somewhat reduced by improving the gas-solid contact and the temperature distribution in the reactor through the better mixing and circulation of reagents. CONCLUSIONS
(i) The ACFB-based gas and power cogeneration process utilizes coal in a simple, effective and environmentally clean manner. (ii) The first- and second-law efficiencies for fuel-gas generation are 35.0 and 27.6%, respectively. The heating value of the gas is 11 MJ/Nm3 (medium); the gas yield is 0.8 Nm3/kg-coal. Desulphurization is achieved by CaS-based sulphur capture during limestone addition to the gasifier bed. (iii) The efficiencies for electric power generation by the standard non-reheat steam cycle are 15.4% (first-law) and 14.6% (second-law). (iv) The overall efficiencies for the gas and power cogeneration are 50.4% (first-law) and 42.2% (second-law) and may be further increased by improvements of design and plant components. REFERENCES
1. P. Nowacki ed., Coal Gusi&ation Processes, Noyes Data Corporation, Park Ridge, NJ ( 1981). 2. L. D. Smoot and P. J. Smith, Coal Combustion and Ga@cation, Plenum Press, New York, NY (1985). 3. P. Basu and S. A. Fraser, Circulating Fluidized Bed Boilers: Design and Operations, Butterworth-Heinemann, Stoneham, MA (1991). 4. X. C. Xu and X. Y. Zhang, Proc. 2nd Int. Symp. Coal Comb., pp. 8-22, China Machine Press, Beijing (1991). 5. K. F. Cen, M. X. Fang, Z. Y. Luo, M. J. Ni, G. Y. Chen, X. T. Li, and B. Y. Yao, Proc. 1st Int. Con$ on CCPG, pp. 240-248, Technical Univ. of Nova Scotia, Halifax (1994). 6. S. V. Makarytchev and K. F. Cen, Proc. 1st Int. ConjI on CCPG, pp. 298-305, Technical Univ. of Nova Scotia, Halifax ( 1994). 7. C. Y. Wen and S. Dutta, in Coal Conversion Technology, C. Y. Yen and E. S. Lee eds., pp. 57-87, AddisonWesley Publishing Company, Reading, MA ( 1979). 8. T. J. Kotas, The Exergy Method of Therm& Plant Analysis, Butterworth, Boston, MA (1985). 9. S. V. Makarytchev, G. D. Smekhov, and M. S. Yalovik, Trans. Russian Acad. Sci., Fluid h4ectumics (in Russian) 1, 155 (1992). 10. S. V. Makarytchev, G. D. Smekhov, and M. S. Yalovik, Trans. Russian Acud. Sci., Fluid Mechanics (in Russian) 6, 157 ( 1992). 11. V. P. Glushko ed., Thermodynamic Properties of Pure Substances (in Russian), Science Press, Moscow ( 1978). 12. S. V. Makarytchev, K. F. Cen, and Z. Y. Luo, Energy-The International Journal 19, 947 (1994). 13. S. V. Makarytchev, K. F. Cen, Z. Y. Luo, and X. T. Li, Chem. Engng. Sci. 50, 1401 (1995) accepted. 14. M. D. Burghardt, Engineering Thermodynamics with Applications, Harper & Row, New York, NY (1982). 15. P. K. Nag, D. Raha, and P. Basu, Proc. 1st Int. Conj on CCPG, pp. 327-344, Technical Univ. of Nova Scotia, Halifax ( 1994). 16. B. Leckner and L. Amand, Proc. 9th Int. ConjI FBC, pp. 891-897, ASME (1987). 17. W. Lin, M. K. Senary, and C. M. Van den Bleek, Proc. I Ith Int. ConjI FBC, pp. 649-654, ASME ( 199 1). NOMENCLATURE
c = Specific heat 8 = Exergy AH = Enthalpy increase HV = Heating value Q = Heat (losses) AQ = Heat increase AT = Temperature difference A W = Turbine work Y = Gas yield Subscripts
C = Carbon coal = Raw coal fg = Flue gas gas = Fuel gas gs = Fuel gas and steam in = Input loss = Heat losses 0 = Overall
power = Electric power s = Solids st = Steam Superscripts
I = First-law value II = Second-law value B = Boiler C = Combustor CG = Combustor-gasifier (difference) EC = Economizer G = Gasifier SH = Superheater ST = Steam turbine Greek letters (Y= Conversion parameter
n = Efficiency
Enrr~’ Vol. 20, No. 12, pp. 1X-1283. 1995 Copyright 0 1995 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0360.5442/95 $9.50 + 0.00
0360-5442(95)00063-l
THERMOCHEMICAL ANALYSIS OF ACFB-BASED POWER COGENERATION S. V. MAKARYTCHEV,?
GAS AND
M. X. FANG, X. T. LI, Z. Y. LUO, and K. F. CEN
Institute of Thermal Power Engineering, Zhejiang University.
3 10027 Hangzhou. People’s Republic of
China (Received 23 December
1994; received for publication 7 June 1995)
Abstract-A first- and second-law analysis is presented for a process developed for simultaneous generation of a fuel gas and electric power (gas and power cogeneration) based on atmospheric circulating-fluidized-bed (ACFB) combustion of coal. The mathematical model has a zone structure, multi-species equilibrium calculations for applicable zone conditions at high temperatures (50 gas-phase and seven solid-phase chemical species) and the concept of freezing of the gas composition at low temperatures. Our analysis shows that the process utilizes coal in a simple, effective and environmentally clean manner. The first- and second-law efficiencies of the process are, respectively, 35.0 and 27.6% for gas generation, 15.4 and 14.6% for power generation, 50.4 and 42.2% overall. The heating value of the gas is 11 MJ/Nm3 (medium). Desulphurization is achieved by using CaS-based sulphur capture during limestone addition to the gasifier bed. Results are compared with data from a 150 kg of coal/h experimental plant.
INTRODUCTION of replacing liquid and gaseous fuels derived from petroleum and NG by synthetic fuels from coal is generally recognized. The technological basis for utilizing coal in combustion and converting it to fuel gas has been created. I,* However, in view of the wide technological gap between industrialized and developing countries, it is unrealistic to expect worldwide use of these highly effective but expensive technologies in the near future. In some rapidly developing countries (e.g., China), there is a need for new gas- and power-generation techniques adapted to the existing technological level. Recently, much effort has been expended in China to develop circulating fluidized bed (CFB) technology to utilize vast national coal resources in an inexpensive, effective and environmentally clean manner. CFB-based power generation in China is now reaching the status of a commercial technology.4 Based on experience in CFB-related research and development, the Institute of Thermal Power Engineering of Zhejiang University started a project aimed at development of a process for simultaneous generation of a fuel gas and electric power (gas and power cogeneration). The first stage is based on ACFB combustion of coal leading later to a pressurized operation. A description of the 150 kg,,,,/h experimental ACFB-based gas/steam generator (the key component of the gas and power cogeneration plant) has been given by Cen et al,’ and a process analysis was carried out by Makarytchev and Cen.‘j A 14 t_,,/h plant is currently under construction in Yangzhong county, Jiangsu province, China, and is scheduled to start operation by the end of 1995. In this paper, we present an analysis of an ACFB-based gas and power cogeneration plant for commercial applications. Because of the heterogeneous nature of the process, the analysis is based on a thermochemical approach, i.e. changes of the thermodynamic quantities within the system (enthalpy, entropy, exergy, etc.) are derived from transformations of chemical compositions and the thermodynamic state of the multicomponent energy carriers, which include fuel and flue gases, solids, and water or steam. The emphasis is on gas parameters and plant efficiencies.
The desirability
PROCESS DESCRIPTION Figure
gas/steam
1 shows the general configuration of a gas and power cogeneration plant. It comprises a generator and a steam turbine cycle. The steam cycle of the plant is a traditional one; it
tvisiting research professor (formerly Moscow University, Russia). To whom all correspondence should be addressed. Present address: Department of Chemical Engineering, University of Sydney, N.S.W. 2006, Australia. 1271
1272
S. V. Makarytchev
et al
Fuel gas Pump
A
I Steam Solids Flue gas to stack N Feed water
Char 1
1
ACFB combustor
Air
Fig. 1. Configuration
ST
Boiled superheater
Steam
of the gas- and power-cogeneration
plant.
consists of a steam turbine, a condenser, a feed pump and a steam rising circuit (economizer, boiler and superheater). A dual-fluidized-bed gas/steam generator (Fig. 2) has two interconnected reactors: an FB gasifierpyrolyzer and an ACFB char combustor. The gasifier bed is fluidized by the mixture of steam, supplied by an internal steam generator, with a recycled part of a fuel gas. The combustor bed is fluidized by air from a compressor. Crushed coal is supplied to the gasifier along with limestone for coal sulphur retention. In the gasifier, coal is heated, pyrolyzes, and reacts with the steam. Hot fuel gas is cooled in an economizer section of a convective pass and, after filtration and dehydration, it is accumulated in storage for future utilization. The char formed during coal pyrolysis is moved to the char combustor by the circulating inert solids and is burned there. The upper part of the combustor contains a superheater for a steam-rising circuit. After the superheater, hot flue gas is separated from solids by a cyclone-type separator and is directed to the convective pass, which contains an internal steam generator and a boiler. The hot solids from the separator recycle to the gasifier and provide heat needed for coal pyrolysis and gasification. The thermal balance of the dual-fluidized-bed combustor-gasifier system may be described as follows. An overall plant thermal input with coal HVcoal is split by the gasifier into two parts: a combustor thermal input fi, (char) and a gasifier thermal input E,, i.e. HVcoal= *” + @” .
(1)
The ratio of fi, to fi, is closely related to the gas-generation process and will be considered below (see Table 3). The heat released in the combustor is used for heating circulating solids AQs, raising steam in the superheater Ae,H, and increasing the flue-gas enthalpy AHfg, viz., H$ = AQs + Ae,n + AHfg .
(2)
I213
ACFB-based gas and power cogeneration ACFB combustor
Gas/solid seperator
Internal steam generator
Flue gas
Fig. 2. Layout of the gas/steam generator.
The gasifier converts its thermal input and the heat delivered by the circulating solids to the fuel-gas heating value HV,,, and a fuel-gas enthalpy AH,,,, i.e. en + AQs = HV,,, -I- AHgaS.
(3)
The combustor-gasifier temperature difference AltG and the solids circulation rate must be sufficiently great to provide the required heat transfer from the combustor to the gasifier. The optimal temperature for the CFB combustor is determined by the level of sulphur retention and an acceptable NO, emission level; it is known to be 850-950°C3 The lower limit of the gasifier temperature is dictated by the rates of pyrolysis and gasification reactions.7 Once the reactor temperatures are chosen, a thermal balance is achieved by adjustment of the solids-to-coal mass-flow ratio, viz.,
kodkgcoa, = AQs/(cs AF=) ,
(4)
where cs is the specific heat of the circulating solids. The present design is based on temperatures of 950°C for the combustor and 8OO’C for the gasifier. For these temperatures and 0.3 mm silica sand with cs = 0.81 kJ/kg K, Eq. (4) yields a lower bound for the solids-to-coal mass-flow ratio of 7.4 kgsorids/kgcoar. The actual ratio should be somewhat higher because of inevitable heat losses. MATHEMATICAL MODEL AND ASSUMPTIONS
The analysis is based on a thermochemical approach, i.e. changes of the chemical compositions and thermodynamic states of the multicomponent energy carriers are used for calculations of the thermodynamic values. The first-law analysis refers to enthalpy transfer within the gas/steam generator and the steam cycle. The second-law analysis relates to exergy transformations.”
1274
S. V. Makarytchev et al
Gas/steam generator
The mathematical model of the gas/steam generator has a zone structure and requires equilibrium calculations for the applicable conditions. In the gasifier and combustor, the chemical compositions of the hot gases are assumed to be in equilibrium at the local temperatures. During the temperature decrease in the heat exchangers of the convective pass, the gas composition changes but remains at equilibrium at the local temperature. At a certain section of the convective pass, the reaction rates become insufficient to achieve equilibrium and the gas composition then remains unchanged downstream. Comparison of the characteristic times for reactions with residence times of the gas shows that chemical composition freezing occurs close to the combustor and gasifier exits. Therefore, the gas composition was treated as frozen starting from the boiler and the economizer sections of the convective pass. The stack temperatures for both fuel and flue gas were taken to be the same at 200°C. The assumed environment for the second-law analysis was as follows: temperature = 25”C, pressure = 1 atm, mole fractions of the gas-phase components are 0.7561 for N1, 0.2034 for 02, 0.0312 for H,O(g), and 0.0003 for CO*. The condensed phase was assumed to be composed of H,O(l), CaC03(s), and CaSO, x 2H,O(s). Calculations for the gas/steam generator were carried out by using computer code for determination of the gas- and condensed-phase species concentrations and thermodynamic functions in multicomponent systems’s” adapted to fossil-fuel applications. For the coal/steam/limestone/air system of concern, the 57 chemical species included were the following: gas-phase Hz, H, CH, CH2, CH3, CH4, C2H, C2H2, C2H3, GH4, C2H5, CzHs, NZ, N, NH, NHZ, NH3, N2H2, CN, COS, NCO, HCN, 02, 0,03, OH, HOZ, Hz02, H20, CO, CO*, HCO, COOH, CH*O, CH30, NO, NO,, NzO, NZ03, N204, NzOs, HNO, SZ, S, SO, SOZ, S03, HS, H,S, Ar; solid-phase C, Ca, CaO, CaS, CaCO,, CaSO, and CaSO,. Data for the species were taken from thermodynamic tables of Glushko. I’Details on chemical equilibrium calculations may be found elsewhere.12*i3 Steam cycle
Assumptions concerning steam-cycle parameters, which are close to design values, are shown in Table 1. Thermodynamic data for steam and water were taken from the steam tables of Keenan et al as given by Burghardt.14
FUEL-GAS GENERATION
Reactions in the gasifier and combustor
When the crushed coal particles are entrained in the steam-blown gasifier, the following reactions occur in succession: coal drying and pyrolysiscoal(s) - msH*O(g ) + Ash(s) + fcC( s) + vmC,H,,OJ,S,( moisture char volatiles
Table 1. Assumptions concerning the steam cycle. Parameter Steam condition pressure ( MPa) temperature (“C) Condenser pressure (kPa) Cooling water temperature (“C) entrance exit Steam turbine isentropic efficiency Feed pump efficiency Heat exchangers efficiency
Value
3.81 450 8.5 25.0 32.0 0.85 0.70 0.80
g);
(5)
ACFB-based
Table 2. Determination
gas and power cogeneration
of the extent of coal conversion Calculated
Experimental composition (wt%)
Gas component
co
I unreacted 26.46 3.49 15.57 0.01 1.88
36.19 5.63 13.56 2.61 2.69 I .69
H? CH., GH, N? I-M other
gas-phase
composition
(wt%) III reacted VM + IO%FC
36.88 8.76 6.25 -
47.13 7.38 10.23 -
1.72
1.60 I .46 16.67 14.26
1.62
0.77 49.66 0.70
31.34 0.23
in the gasilier.
II reacted only VM
1.46
co2
Hz0
VM + steam
1275
19.21 24.93 0.63
1.27
reactions of volatiles with steam-
C,.H,O,N,S, heterogeneous
+ (c-fc-o)HIO
carbon-steam
-
(c-fc)CO
+ (c-fc+h/2-o-s/2)H,
+ (n/2)N,
+ sH,S;
(6)
reactions (char gasification)-
C(s) + I-VW - CO(g) + Hz(g) CO(g) + H,O(g) us> + CO*(g) -
(7)
9
CO, + H,(g) 9
(8)
2CWg) .
(9)
In the preceding equations, C(s) is solid carbon bound in the char along with ash. The coefficients ms, fi, vm, and c, h, o, n, s in Eqs. (5) and (6) denote, respectively, the mole fractions of moisture MS, fixed carbon FC, volatile matter VM, and individual elements C, H, 0, N, and S in the raw coal. These coefficients are determined by the coal composition (see Table 3). The volatiles are composed of gaseous CH4, C2H6, CO, CO*, HZ, H20, H2S, NH,, and some other compounds. Because of the short residence time, the slow heterogeneous reactions (7)-(9) are not completed in the gasifier. Along with the circulating inert solids, the unreacted char moves to the combustor where heterogeneous char combustion occurs as follows: Table 3. Fuel characterization Reactor Parameter
Plant
Gasifier
Combustor
Fuel type Fuel composition
raw coal
VM+MS
FC + Ash
50.56 25. I 1 2.71 21.62 63.39 3.88 6.51 1.13 0.76 24.58 1.00 24.58 -
-
70.05
(wt%) FC VM MS Ash C H 0 N S Heating value (MJ/kg,.,)
Mass input (kg,,&,,,) Thermal input (MJ/kg,,,,,) Theor. air (kg,,$kg,,,) (kg,i,lkg,,,) Theor. steam (kg.,,.,/kg,.,) (kg,,,.,/kg&
90.26 9.74 46.12 13.95 23.40 4.06 2.73 28.75 0.278 8.00 0.43 0.12
-
29.95 70.05 -
22.91 0.722 16.58 8.06 5.82 -
S. V. Makqtchev
1276
C(s)
+ O*(g)
C(s) + (wwg) CO(g) + (1~2Mg)
et al
-
CO,(g)
(10)
9
+ CO(g) 9
(11)
-
(12)
a(g)
*
Extent of coal-to-fuel-gas conversion The extent of this conversion and the chemical composition of the fuel gas depend on the gasifier temperature, pressure and the residence times of reactants. For heterogeneous gas-solid reactions, additional important factors are coal reactivity, heating rate, reactive surfaces, and surface temperatures. The design operating temperature of the gasifier (NWC) and the residence time of the coal particles (approximately, 5 min) are sufficient for complete pyrolysis [reaction (5)] but not for completion of the char-gasification reactions (7)-(9), which are slow at 8OO’C.’ To determine the extent of reaction in the gasifier, the composition of the fuel gas generated in a 150 kg,,,,/h experimental gas/steam generato? was compared with equilibrium calculations. Table 2 shows the experimentally observed gas composition and the composition calculated for the three suc-
cds rslio 0
t
i 10’
0
0.4
0.8
1.2
1.6
2.0
Gasifier steam ratio 1.0
(b)
3 E g
0.5
2
Cold 9~ efficisacy
.i 3
Gasifies: MOT! Combttstor 95OT. air rstio 1.2
0
-
I
I
I
I
0.4
0.8
1.2
1.6
20
2.0°
Gasifier steam ratio Fig. 3. Effect of gasifier steam on fuel-gas composition (a) and coal-to-gas conversion parameters(b).
1277
ACFB-based gas and power cogeneration
cessive gasification steps: (I) only pyrolysis reaction (5) is completed (unreacted VM and steam); (II) the gas-phase reaction (6) of the volatiles with steam is completed (only VM is reacted); (III) the gassolid char-gasification reactions (7)-(9) are partially completed with 10% of the total char mass gasified (VM and lO%FC are reacted). The assumption for step III was used previously by Nag et alI5 for analysis of a pressurized, air-blown, Lurgi-type FB gasifier. Table 2 indicates that for atmosphericpressure, steam-blown gasification at 800°C, the observed concentrations approach the calculated values for step II. This means that the gas-phase reactions are close to completion while the extent of char gasification is negligible. The gasifier acts as a pyrolyzer-carbonizer. In view of these results, we have based the plant analysis on the following assumption: the fuel gas at the gasifier exit is an equilibrium volatiles-steam mixture at 800°C after completion of reactions (5) and (6) and all of the char formed during pyrolysis is burned in the combustor. For these assumptions, fuel characteristics for the gasifier and combustor are given in Table 3. Theoretical requirements for gasifier steam and combustor air are determined according to reactions (6) and (lO)-( 12), respectively. Fuel-gas
parameters
The main parameters determining gasifier performance are the fuel-gas composition, the contents of carbon in the fuel gas (kg/Nm3), the gas heating value HV,, (kJ/Nm3), and the gas yield YgaS (Nm’/kg,,,,). The main dimensionless parameters are the carbon conversion (ac) and the cold fuelgas efficiency niaS(first-law value); these are defined as follows: ffc = carbon-in-fuel-gas x Y&carbon-in-coal ,
(13) (14)
rltas = HV,,S Y&?SlHV,,,. Using the data of Table 3, carbon conversion in the gasifier is approximated as oc = (C-FC)/C = (63.39-50.56)/63.39 = 0.20 .
(15)
The fuel-gas composition and cold fuel-gas efficiency as functions of the gasifier steam ratio (actual steam/stoichiometric steam) are shown in Figs. 3(a) and 3(b). With an increase of the steam ratio, the concentrations of the combustible species CO and CH, in the fuel gas decrease, while those of the oxidation products HZ0 and CO* rise [Fig. 3(a)]. This explains a decrease in the gas heating value from 16 MJ/Nm3 for zero steam ratio (dry coal pyrolysis) to 9 MJ/Nm3 for a steam ratio of 2.0 [Fig. 3(b)]. At the same time, the gas yield increases from 0.5 to Sulphur capture --------ce __ ___--_-
106
,
.@
1.0
#*
s c ,a 0.4 g ?
-
‘J
stoam rrlio 1.O
0
1
2
3
4
Co/S molar trtio Fig. 4. Effect of limestone on H2S content in the fuel gas.
5
1278
S. V. Makarytchev et
al
1.ONm-‘/kg,,,,. As the result, & remains almost unchanged at 35% for the steam-ratio range O-2.0. At the design steam ratio of 1.O,IN,,, = 10.9 MJ/Nm3 (2600 kcal/Nm”), which corresponds to medium heating-value gases and agrees well with the value 10-12 MJ/Nm3 for the experimentally generated gas.5 Desulphurization
During pyrolysis, volatile sulphur compounds are produced as H,S and a small amount of COS. Even for low-sulphur coal, the amount of H2S in the fuel gas is about 2% [Fig. 3(a)], which poses an environmental issue. Desulphurization of the fuel gas is achieved by adding a Ca-based sorbent (in the present design, limestone) directly to the gasifier bed. At the gasifier temperature, limestone calcifies, i.e.
0.8
0.2
0
0.4
0.8 Gasifier
1.2
1.6
2.0°
steam ratio
(b) 1.0
0.8
0.6
u.6
1.0
1.2
I .4
1.6”
Combustor air ratio Fig. 5. Effects of gasifier steam (a) and combustor air (b) on the thermal balance of the gas/steam generator.
ACFB-based
gas and power cogeneration
1279
CaO( s) + COz( g) .
(16)
CaCO,( s) CaO then reacts with H2S as follows: H&g)
+ CaO( s) -
H,S(g) + (3/2)0,(g)
CaS(s) + H,O(g)
,
(17)
SO,(g) + H,O(g)
7
(18)
-
SO,(g) + CaO( s) + 3CO( g) -
CaS( s) + 3C02( g)
(19)
In an oxygen-deficient gasifier atmosphere, the amount of SO2 is negligible and sulphur is primarily converted to solid CaS. Sorbent particles containing CaS are transferred by the circulating solids to the combustor where CaS is oxidized to solid CaSO, by the process
Condensate 4.1
I
Combustor
Economizer
v////////-//m
Fuel gas
I
Boiler + superbeater
I
Steam turbine
-
I
I
Power 15.4 Fig. 6. Plant heat-flow diagram.
Interoal
Stackgas and loss
S. V. Makarytchev et al
CaS( s) + 202(g) - CaSO,( s)
(20)
and are then disposed of with the ash. Figure 4 shows the calculated amounts of H$, CaO, CaS in the gasifier and the fractional sulphur capture as CaS vs Ca/S molar ratio of an added sorbent. By increasing the amount of sorbent, the H,S concentration may be reduced to meet environmental constraints. However, increased sorbent addition results in larger amounts of solid waste and also leads to an increase in NO, emissions.‘6*‘7The optimal value for the Ca/S ratio is in the range 2-2.5. Details of the CaS-based desulphurization in a fluidizedbed reactor are considered elsewhere.13 RESULTS OF THE FIRST-LAW
ANALYSIS
Thermal petiormance of the gas/steam generator
The gas/steam generator converts the plant thermal input HVcoalto the fuel-gas heating value HV,,, in the gasifier and to steam heat in the economizer AHzc, boiler AZ,, and superheater AHzH.The thermal balance equation for the gas/steam generator is HV,,,, = HV,,, + W?
+ A@, + A@? + Q,oss9
(21)
where Qlossis the total heat loss in the heat exchangers and with the stack gases. Figures 5(a) and 5(b) show the normalized on HV,,, terms of Eq. (21) as functions of a gasifier steam ratio [Fig. 5(a)] and a combustor air ratio [Fig. 5(b)]. The relation between the terms of Eq. (21) is stable over a wide range of operating parameters. The increase of Qlossat the air ratio less than unity is caused by the rise of the content of combustible matter (mainly CO) in the flue gas. In a design operating regime (gasifier: 8OO”C,steam ratio 1.0; combustor: 95O”C, air ratio 1.2), 35% of the HVcoa, is converted to the HV,,. 45% is used for rising steam (2, 17 and 26% in economizer, boiler and superheater, respectively), and 20% is lost in the heat exchangers and with the stack gases. The first-law thermal efficiency of the gas/steam generator T& is the sum of the gas-generation (vi,,) and steam-generation efficiency ($), viz.,
it equals 0.35 + 0.45 = 0.80. 80 -
- 80 Comburtor: 9SO°C Guifier: 800°C, rtoam mtio 1.0
60-
-60
2 6 I 3 s u 1
Ovomll officioncy
40 -
-
Gas generation
Powor gonoration
0 0.6
I
I
I
I
0.8
1.0
1.2
1.4
Comburtor
air ratio
Fig. 7. The effect of combustor air on the plant thermal efficiency.
-0 1.6
40
ACFB-based
1281
gas and power cogeneration
Gasifier 1
I
Condc
r
Solids
oadenser 1.7 Power 14.6
Fuel gas
Fig. 8. Plant exergy-flow
Power generation
diagram.
and overall plant eficiency
Figure 6 illustrates the energy conversion process within the plant. The numbers in Fig. 6 correspond to the enthalpy values of the different energy carriers in the design operating regime. Regarding the steam cycle efficiency, only 31% of the superheated steam heat ( 15.4% of the HV,,,,) is converted to the steam turbine shaft work AWsTwhile the rest is withdrawn by the cooling water in the condenser. The first-law-based indexes of plant performance are the cold fuel-gas efficiency qk,, [Eq. (14)], power-generation efficiency
and overall plant efficiency I I I qo = qgas + qpower ’
(24)
The values of these efficiencies are almost independent of the gasifier steam ratio. Their dependence
S. V.
1282
Makarytchev et al
on combustor air ratio is shown in Fig. 7. In the design regime, the first-law efficiencies are 0.350 for gas generation, 0.154 for power generation, and 0.504 overall. RESULTS OF THE SECOND-LAW ANALYSIS
Exergy conversion,
losses and irreversibilities
Figure 8 shows the process of the exergy conversion in the gas and power cogeneration plant under the same conditions as for Fig. 6. Black triangles in Fig. 8 indicate the destruction of the exergy due to irreversibilities in the plant components. Table 4 gives the summary of the exergy analysis. Results show that 27.6% of the coal chemical exergy Scoalconverts to chemical exergy of the fuel gas ggas and 14.6% to steam turbine work AWsT. The exergy losses through the effluent streams are insignificant: 3.4% with the stack gases and 0.3% with the cooling water. This is in contrast with the corresponding first-law-based data: 5.8 and 24.0% of the HV,,,, respectively. The main source of the exergy loss (50% of the ‘%,,,,) are the irreversibilities within the gas/steam generator. Two-thirds of this amount is the result of imperfect heat transfer in the heat exchangers, and the rest is an internal irreversibility of the gasification and combustion processes. Second-law eficiency for gas and power cogeneration
Using data of Table 4, the second-law efficiencies may be evaluated as follows: qi= = C%gaslc&coa, = 0.276 , II
qPwer = AWsTIZ,,, = 0.146 I$
=
qias
+
q;,,,
=
(25) ,
(26)
0.422 .
(27)
Because of irreversibilities, these efficiencies are smaller than the corresponding first-law efficiencies. To improve process efficiency, there must be a decrease in the irreversibility of gas-to-steam heat transfer in the boiler and the superheater (32% in the total exergy balance). This may be accomplished by avoiding excessive heat dissipation into the environment by using better isolation materials and by appropriate matching of the temperatures of heating and heated streams. The irreversibilities in gasifier (7.5%) and combustor (9.8%) are inevitable features of the related chemical transformations. They can
Table 4. Summary of the exergy analysis. Parameter Exergy input with coal
26007
100.0
Exergy conversion fuel gas chemical exergy turbine work output total conversion
7174 3797 10971
27.6 14.6 42.2
892 77 969
3.4 0.3 3.7
1948 2524 5447 2875 252 669 338 44 14067
7.5 9.8 20.9 10.9 0.9 2.6 I.3 0.2 54. I
Exergy losses with the stack gases with the cooling water total losses Irreversibilities gasifier + economizer combustor superheater boiler internal steam generator turbine condenser feed pump total irreversibil.
ACFB-based gas and power cogeneration
1283
be somewhat reduced by improving the gas-solid contact and the temperature distribution in the reactor through the better mixing and circulation of reagents. CONCLUSIONS
(i) The ACFB-based gas and power cogeneration process utilizes coal in a simple, effective and environmentally clean manner. (ii) The first- and second-law efficiencies for fuel-gas generation are 35.0 and 27.6%, respectively. The heating value of the gas is 11 MJ/Nm3 (medium); the gas yield is 0.8 Nm3/kg-coal. Desulphurization is achieved by CaS-based sulphur capture during limestone addition to the gasifier bed. (iii) The efficiencies for electric power generation by the standard non-reheat steam cycle are 15.4% (first-law) and 14.6% (second-law). (iv) The overall efficiencies for the gas and power cogeneration are 50.4% (first-law) and 42.2% (second-law) and may be further increased by improvements of design and plant components. REFERENCES
1. P. Nowacki ed., Coal Gusi&ation Processes, Noyes Data Corporation, Park Ridge, NJ ( 1981). 2. L. D. Smoot and P. J. Smith, Coal Combustion and Ga@cation, Plenum Press, New York, NY (1985). 3. P. Basu and S. A. Fraser, Circulating Fluidized Bed Boilers: Design and Operations, Butterworth-Heinemann, Stoneham, MA (1991). 4. X. C. Xu and X. Y. Zhang, Proc. 2nd Int. Symp. Coal Comb., pp. 8-22, China Machine Press, Beijing (1991). 5. K. F. Cen, M. X. Fang, Z. Y. Luo, M. J. Ni, G. Y. Chen, X. T. Li, and B. Y. Yao, Proc. 1st Int. Con$ on CCPG, pp. 240-248, Technical Univ. of Nova Scotia, Halifax (1994). 6. S. V. Makarytchev and K. F. Cen, Proc. 1st Int. ConjI on CCPG, pp. 298-305, Technical Univ. of Nova Scotia, Halifax ( 1994). 7. C. Y. Wen and S. Dutta, in Coal Conversion Technology, C. Y. Yen and E. S. Lee eds., pp. 57-87, AddisonWesley Publishing Company, Reading, MA ( 1979). 8. T. J. Kotas, The Exergy Method of Therm& Plant Analysis, Butterworth, Boston, MA (1985). 9. S. V. Makarytchev, G. D. Smekhov, and M. S. Yalovik, Trans. Russian Acad. Sci., Fluid h4ectumics (in Russian) 1, 155 (1992). 10. S. V. Makarytchev, G. D. Smekhov, and M. S. Yalovik, Trans. Russian Acud. Sci., Fluid Mechanics (in Russian) 6, 157 ( 1992). 11. V. P. Glushko ed., Thermodynamic Properties of Pure Substances (in Russian), Science Press, Moscow ( 1978). 12. S. V. Makarytchev, K. F. Cen, and Z. Y. Luo, Energy-The International Journal 19, 947 (1994). 13. S. V. Makarytchev, K. F. Cen, Z. Y. Luo, and X. T. Li, Chem. Engng. Sci. 50, 1401 (1995) accepted. 14. M. D. Burghardt, Engineering Thermodynamics with Applications, Harper & Row, New York, NY (1982). 15. P. K. Nag, D. Raha, and P. Basu, Proc. 1st Int. Conj on CCPG, pp. 327-344, Technical Univ. of Nova Scotia, Halifax ( 1994). 16. B. Leckner and L. Amand, Proc. 9th Int. ConjI FBC, pp. 891-897, ASME (1987). 17. W. Lin, M. K. Senary, and C. M. Van den Bleek, Proc. I Ith Int. ConjI FBC, pp. 649-654, ASME ( 199 1). NOMENCLATURE
c = Specific heat 8 = Exergy AH = Enthalpy increase HV = Heating value Q = Heat (losses) AQ = Heat increase AT = Temperature difference A W = Turbine work Y = Gas yield Subscripts
C = Carbon coal = Raw coal fg = Flue gas gas = Fuel gas gs = Fuel gas and steam in = Input loss = Heat losses 0 = Overall
power = Electric power s = Solids st = Steam Superscripts
I = First-law value II = Second-law value B = Boiler C = Combustor CG = Combustor-gasifier (difference) EC = Economizer G = Gasifier SH = Superheater ST = Steam turbine Greek letters (Y= Conversion parameter
n = Efficiency